The present invention relates to solenoids. More particularly, it relates to solenoids whose pole faces are shaped for controlled force characteristics. Most particularly, it relates to solenoids for direct actuation of automotive valves, achieving efficient pull-in from a distance, rapid deceleration on approach to closure, and a rapid increase to a latching force, with controllable electromechanical damping associated with latching and unlatching.
The following discussion, and the invention to be described, relate particularly to solenoids that close magnetically with high speed, high magnetic force, and a large area of mating magnetic pole faces. Solenoids generating large magnetic forces are especially prone to close with an impact, generating noise and damage to the solenoid. The motions in solenoids and magnetic bearings have been successfully controlled by servo feedback control under limited conditions. For gradual motions and where the magnetic gap does not change by large fractional amounts, stable servo control has been demonstrated. When the magnetic gap changes from open to closed with mechanical poleface contact, and when the motion toward closure is very rapid, servo control is very difficult to achieve, especially for obtaining quick closure to a low-impact landing. This difficult control problem needs a solution if there are to be commercially viable automotive electric valve actuation systems, for example.
The invention to be disclosed in the following specification addresses the solenoid control problem not by way of a new servo controller, but by a fundamental redesign of the solenoid itself. The broad objective of the invention is to achieve electromagnetic characteristics that simplify the task of servo control. To understand the criteria behind this solenoid redesign, it is necessary to gain some understanding of the electrical servo control problem itself. The remainder of this background section examines the control problems that, in the prior art, have been addressed almost entirely from the controller side. This examination provides more detail than would normally be devoted to prior art, including issues that may not be widely understood. The intent of this close examination is to define the context of the current invention in sufficient detail that the design criteria will be understood.
A difficulty with servo control in a magnetic solenoid is that, in order for a controller to change electromagnetic force rapidly, the entire energy residing in the solenoid magnetic field must be altered in roughly the same proportion that the force is altered. To close a solenoid, electrical energy applied to windings is converted to magnetic field energy, which can be considered to reside mostly in the air gaps between the ferromagnetic armature and the magnetic yoke, which is part of the stator structure. For a given flux density, the magnetic energy in high-permeability core material is much lower than in air, in proportion to the relative permeability of the core material (typically in excess of 1000). If the total magnetic flux linking the yoke windings is held constant, then the magnetic field energy will diminish as the armature-to-yoke gap closes, eliminating the high-energy field passing through air. In geometries where armature motion is predominantly parallel to the magnetic flux crossing an air gap, the magnetic force of attraction between the armature and yoke is approximately the partial derivative of magnetic field energy with respect to armature position, evaluated under the constraint that total flux is held constant. As a flux-carrying solenoid gap closes, one can say that the field energy in the gap is transformed into mechanical energy. Conversely, when a narrow flux-carrying solenoid gap is mechanically forced to open, mechanical energy is transformed into magnetic field energy as flux-filled magnetic gap volume is created. In the idealized case of zero applied winding voltage and zero winding resistance, e.g., the case of a shorted superconductive solenoid winding with no source or sink of electrical energy, then pulling an armature away from a magnetized yoke will result in a growing armature-yoke gap filled with magnetic energy, that energy coming entirely from the mechanical work that pulls the armature. A difficulty with servo control of solenoid motion through control of magnetic force is that, in order to change force rapidly, a reservoir of magnetic energy must be filled or depleted rapidly. The magnitude of the energy in that reservoir is on the order of the total energy that would be delivered to the armature if the magnetic gap were closed completely, starting from the gap width for which the energy transfer problem is defined.
While the perspective of energy conversion sheds light on the bounds of possibility in a problem of this sort, viewing the problem in terms of momentum is sometimes more helpful for the control details. xe2x80x9cMagnetic momentumxe2x80x9d may be understood by direct analogy to mechanical momentum. The expression for mechanical kinetic energy, xc2xd MV2 for mass M and velocity V, has its magnetic energy counterpart in xc2xd LI2 for inductance L analogous to mass and current I analogous to velocity. Differentiating energy with respect to velocity or its analog, current, yields mechanical and magnetic momentum expressions: MV for mechanical momentum, and LI for magnetic momentum. Inductance L expresses the electrical xe2x80x9cinertiaxe2x80x9d that resists change in current I, and when this inertia is overcome by the continued application of voltage, magnetic momentum LI is altered. A basic equation states that xe2x80x9cLI=n"PHgr",xe2x80x9d where xe2x80x9cnxe2x80x9d is the winding turns count and xe2x80x9c"PHgr"xe2x80x9d is the total magnetic flux linking the n windings. It is common to refer to the product xe2x80x9cn"PHgr"xe2x80x9d as the flux linkage. One sees that xe2x80x9cflux linkagexe2x80x9d is an alternative expression for xe2x80x9cmagnetic momentumxe2x80x9dxe2x80x94the different terms emphasize different aspects of the same quantity. There is thus an xe2x80x9cinertiaxe2x80x9d associated with magnetic flux, making flux difficult to change rapidly. The scaling of this magnetic inertia depends on the winding count, n. By extension of the mechanical analogy, a source of mechanical force and motion may transfer energy to a moving mass with a selectable mechanical advantage or disadvantage, established by a fulcrum or gear ratio. The mechanical advantage is selected for an impedance match between the characteristics of the energy source and the load. In a solenoid, by analogy to a mechanical gear ratio, the winding count n establishes the xe2x80x9celectrical advantagexe2x80x9d exerted by an electronic driver on the magnetic system. If n is low, the electronic driver has a high xe2x80x9cadvantagexe2x80x9d analogous to xe2x80x9clow gearxe2x80x9d in a mechanical transmission, so that the driver can alter magnetic flux rapidly for a given voltage output. The problem associated with a low winding count is a high current draw, analogous to a motor having to rev at high speed when driving a load in low gear. Raising the winding count reduces the current draw for a given combination of magnetic flux and field gap, and it also reduces the ability of the driver to alter magnetic flux rapidly. As will be shown, an objective of the present invention is to cause needed changes in magnetic force, in order to direct the course of a solenoid armature toward soft landing, without demanding rapid changes in the total magnetic flux linkage, n"PHgr". With a reduced tendency toward rapid xe2x80x9cflux slewing,xe2x80x9d i.e. ramping flux up or down at a rate limited by power supply voltage, the electronic controller can gain better control of mechanical motion. Some of this control advantage can be traded off for a reduced peak current draw, by increasing the winding count, n. The improved electromagnetic design does not xe2x80x9csolvexe2x80x9d the control problem, but it makes it far more tractable, within engineering and economic constraints.
In order to control solenoid force and therefore cumulatively influence mechanical momentum, one must cumulatively vary magnetic momentum in order to change force. One sees that an extra xe2x80x9clayerxe2x80x9d of inertia and energy accumulation is inserted into the control path from actuation voltage to controlled position, resulting in a third-order control system. Consider a fast-moving solenoid with a high transfer of energy per stroke, as exemplified by either side of a double-acting automotive valve solenoid. FIG. 1 illustrates such a solenoid, typical of the prior art, in end elevation section. A powerful spring system, not shown, restores the ferromagnetic armature 120 to the midpoint between two latching yokes, 100 and 180. Top yoke 100 consists of an E-core 105 of ferromagnetic material, typically stacked laminations lying flat in the plane of the diagram, plus a winding 110 wrapped around the center post of the xe2x80x9cUxe2x80x9d and seen in section on the left and the right of the center post. Magnetic flux passing through air is concentrated predominantly across the gaps between the armature and the yoke posts, going in one direction (e.g., up) via the center post at 111 and going in the opposite direction (e.g., down) via the side posts 116 and 117 on the left and right. When 120 is pulled off center and latches against 100, gaps 111, 116, and 117 close, while pulling 120 in the opposite direction and latching against yoke 180 causes gaps 111, 116, and 117 to double in width compared to the FIG. 1 illustration. FIG. 2 illustrates a very similar solenoid in a U-core topology. Here, yoke 200 consists of stator 205 and a pair of windings, 210 and 215, wrapping around the right and left posts of the U-core. Windings 210 and 215 would typically be wound together, in series or in parallel, to connect to a single pair of terminals. Yoke 280, opposite armature 220 from yoke 200, is similar to yoke 200 and provides for latching 220 on the far side from 200, as with yoke 180 in relation to yoke 100. The U-core topology with armature 220 has two symmetric air gaps, 211 and 216, from the left and right posts of the U-core. In FIG. 3, the armature 220 has been pulled up and latched against U-core yoke 200, causing open air gaps 211 and 216 to become closed air gaps 311 and 316.
The following operational description for the armature and yoke assembly of FIGS. 2 and 3 is applicable with little change to the armature and yoke assembly of FIG. 1. More subtle differences in space utilization, latching force, and efficiency exerting a force on the armature across a substantial magnetic gap, will be discussed later. When the armature 220 is magnetically latched on either side, a typical mechanical energy stored in the spring (not shown here, but actually coupled between the armature and static structures holding the yoke) is 1.4 joules. Suppose that armature 220 is latched to top yoke 200, as shown in FIG. 3. Because the magnetic flux goes through closed gaps at 311 and 316, corresponding to the open gaps 211 and 216, without bridging through air, the magnetic energy is low, the inductance is high, and the electric current needed to maintain a latching flux is lowxe2x80x94a property well recognized in the art, where power to maintain latching must be kept low. If the small voltage needed to maintain the latching field is removed, the field will decay relatively slowly, since the current-times-resistance product produces a small voltage to collapse the field. A reverse xe2x80x9cbrakingxe2x80x9d voltage will speed the reduction of field strength, with relatively low power transfer due to the low current. When the magnetic force falls below the applied mechanical forces, including spring force and other load forces (e.g., of gas pressure differentials acting on an automotive cylinder valve), then the armature starts to release, whereupon a growing magnetic gap will initially cause winding current flow to rise, typically even in the presence of an external braking voltage. The spring energy is going partly to impart kinetic energy to the armature and partly to build up magnetic field energy, a portion of which may be xe2x80x9charvestedxe2x80x9d from windings 210 and 215 using appropriate circuitry. If the magnetic flux is reduced slowly as the armature pulls away, magnetic attraction will persist and oppose much of the spring force that would otherwise be imparting kinetic energy to 220. Induced winding current will also climb quite high if flux is reduced slowly, as electrical energy is dissipated in winding resistance and, possibly, partially harvested by circuitry connected to the windings.
A strong braking voltage will reduce the field strength rapidly, giving a quick release and minimizing transfer of energy from mechanical to magnetic form. If the solenoid release allows an automotive exhaust valve to open against considerable residual cylinder pressure, a substantial fraction, as much as half or more of a typical 1.4 joule spring energy, is dissipated as the valve pushes upstream against outrushing exhaust gases. To get the armature to the opposite side and latched against spring force, the windings must be energized early, e.g., starting even before the armature is halfway across from one yoke to the other. To exert a high force and replace lost energy, e.g., 0.7 joules, in a period of a millisecond or so, the implied average power level works out to 0.7 joules divided by 0.001 seconds, or 700 watts. A higher peak power capability is typically needed. The number of windings chosen for 210 and 215 around core 205 of yoke 100, and similarly for windings in yoke assembly 280, needs to be set so that one to two kilowatts of power can be transferred at a reasonable combination of amperes and volts, e.g., 40 volts and 25 to 50 amps for 1000 to 2000 watts. The peak voltage will set the peak rate of change of the flux linkage n"PHgr" or magnetic momentum, i.e. d(n"PHgr")/dt)=40 volts, approximately, neglecting electrical resistance. If power transfer is extended over one millisecond to transfer 0.7 joules in our example, then the net change in flux linkage n"PHgr" must be the product (d(n"PHgr")/dt)xc2x7(0.001 seconds)=0.04 volt-seconds. A volt-second is a unit of electrical impulse to change magnetic momentum, analogous to a newton-second of force impulse to change mechanical momentum. This volt-second figure to drive a given change in flux linkage remains about the same (discounting resistive voltages) whether flux is being built up for a large gap, to cause a solenoid to close after a substantial energy loss, or for a small gap, to cause a solenoid to latch promptly after a period of deceleration toward a stop. Suppose that peak wattage is to be constrained, to hold down the cost of electrical driver semiconductors. To replenish a given amount of lost valve energy, wattage is held down by using a relatively long time period to transfer a given amount of energy, e.g., one full millisecond for a total armature transit time of 3 to 3.5 milliseconds, as opposed to a much shorter period like 0.2 milliseconds. If the power transfer time is pushed much shorter, the designer encounters limits of magnetic saturation. If the power transfer time is pushed much longer, then yoke 280 would have to start pulling on armature 220 immediately upon the release of 220 from a latched state against yoke 200. Because of problems with flux leakage and spreading of the magnetic field and core saturation limits, it is not feasible to achieve a high force of magnetic attraction across such a large gap. The designer is typically constrained to make up energy losses in a time period somewhat, but not too much, under one-half the transit time of the armature from one yoke to the other.
Continuing with reference to a solenoid operated in conjunction with an automotive valve, as the armature approaches landing when engine RPMs are high, it is often desirable to have landing take place promptly, especially when a valve makes a quick transition from open to full-closed. (When a valve is approaching full-open, it is relatively unimportant how much time is required to get the armature through the last 10% of its travel to a fully latched position.) A transition path for an opening exhaust valve is illustrated by the graph of FIG. 4. 490 indicates the horizontal time axis, labeled xe2x80x9ctxe2x80x9d, for a multi-trace graph of events associated with solenoid opening. The vertical scale on axis 400 is marked in increments from xe2x88x921 to +1. The extreme values of xe2x88x921 and +1 represent the range from xe2x88x92100% to +100% of full scale for three traces: xe2x88x92100% to +100% of center-to-peak travel for trace 420; and xe2x88x92100% to +100% of peak applied voltage for traces 425 and 430. Trace 410 represents exhaust pressure exerted against the valve. Trace 420 represents exhaust valve position, moving from fully closed on the top left to fully open on the bottom right. The first movement of trace 420 begins shortly before vertical wedge-shaped time marker 491. Observe that exhaust pressure trace 410 is falling before trace 420 indicates initial valve opening. In fact, trace 410 extends upward beyond the left side of the time range of FIG. 4 to a much higher peak pressure soon after ignition. The pressure falls as the piston descends, allowing the exhaust gas to expand. If the exhaust valve had not opened, the decline in exhaust pressure differential across the valve would have slowed to a stop at a positive pressure differential and then started to climb again, during the time range of FIG. 4, as the piston reached bottom dead center and started back up. In the events actually depicted in FIG. 4, the opening of the valve causes the rate of decrease in exhaust pressure to accelerate and then slow again as the cylinder pressure approaches equilibrium with exhaust manifold pressure, i.e. zero pressure differential exerting force on the valve. The exhaust pressure at time marker 491 is approximately 8 atmospheres in the illustrated simulation. Trace 425 shows a negative voltage pulse applied to the coil on the top holding yoke, e.g., yoke 200 of FIGS. 2 and 3, with the +1 armature position being the position of armature 220 in FIG. 3, latched to yoke 200 at a displacement of about 4 millimeters above its neutral center position. Trace 435 represents current in the coils associated with the top holding yoke, declining from a low holding value toward zero, the value reached at time marker 491 and where voltage trace 425 switches from its negative extreme to zero. The valve does not begin a rapid opening acceleration until current trace 435 is quite low, both because little current is required to maintain an established magnetic flux for the closed armature-yoke gap, and because the considerable exhaust pressure represented by trace 410 is neutralizing a substantial fraction of the mechanical spring force that would otherwise be pulling the solenoid and valve strongly to open. Voltage trace 425 continues at zero during a quiescent electrical period from time marker 491 to marker 492, at which point trace 425 is replaced by voltage trace 430, the voltage drive on the bottom yoke, e.g., yoke 280 of FIGS. 2 and 3, the yoke that pulls the armature to a magnetically closed bottom position with the valve fully open. Current trace 435 for the upper winding is replaced by trace 440 for the lower winding, with trace 440 starting up from zero at time marker 492. Observe that the application of voltage and the initial ramping up of current begins before the armature has traveled halfway from top to bottom position. Electromagnetic efficiency is very low up to time marker 493, where current trace 440 reaches its maximum, since the solenoid magnetic gap is large. Dashed trace 450, representing flux linkage in the lower yoke, approaches a level of significant core saturation on approach to time marker 493, and the resulting decrease in inductance causes current trace 440 to stop curving down and begin rising more sharply at it approaches a maximum value. A clamp in the controller recognizes a nominal saturation limit for signal trace 450 at +1 on axis 400 and switches to a mode of operation that clamps the flux, whereby drive voltage 430 is reduced to prevent 450 from exceeding the preset limit. When 450 becomes clamped, current trace 440 begins to fall, following the curve that maintains constant flux linkage as the magnetic gap closes and trace 420 comes closer to the full-closed level of xe2x88x921 on vertical scale 400. At time marker 494, the flux linkage clamp logic releases as the motion control logic takes over and calls for a reduction in flux linkage trace 450. This flux reduction is driven by a negative voltage drive on trace 430 from time marker 494 to marker 495. The fall of current trace 440 becomes steeper in this time interval, driven down both by the decreasing magnetic gap and by the decreasing flux linkage. In the interval from time marker 494 to 495, magnetic pull-in force is being reduced to allow the armature to decelerate toward a landing at low velocity. At time marker 495, just before the first landing contact and slight bounce of the armature, the drive voltage kicks back up to the maximum limit, driving flux linkage trace 450 quickly up not quite to saturation, in order to provide enough magnetic force at closure to latch and hold the armature against full spring force. The magnetic gap closes fully, with a slight bounce, about midway between time markers 495 and 496, and at 496, the control system switches to a holding mode, reducing the drive voltage in order to stabilize flux linkage at a holding value. The corresponding holding current of current trace 440 is slightly higher than the initial value of trace 435, which was able to be reduced slightly because exhaust pressure was helping to hold the valve closed, whereas no corresponding pressure is helping to hold the valve open at the right end of trace 440.
Two related electromagnetic design issues are illustrated by the traces of FIG. 4. First, where a substantial fraction of the spring energy launching the armature of an exhaust valve is dissipated in overcoming an opposing flow of exhaust gas, the armature would fall far short of closure on the valve-open side unless a large power pulse were applied to the lower yoke when the magnetic gap was still quite large. The worst-case valve energy losses are overcome with worst-case low electromagnetic efficiency, resulting in high peak power requirements and high net energy drain, much of the energy being dissipated by electrical resistance. Magnetic pulling efficiency at large magnetic gaps is improved if the areas of attracting pole faces can be increased and if the narrowest pole face dimension can be increased. The efficiency loss issues for a given yoke winding are pole face area, spreading of the magnetic field between the pole faces and the armature with large gaps, and shunting of magnetic flux between poles of the yoke without reaching across to the armature. As will be discussed, U-core topologies have advantages over E-core topologies both in reducing flux shunting losses and in reducing the problem of spreading magnetic fields, since the two narrow end poles of an E-core are replaced by a single, wider end pole of a U-core. Thus, U-core topologies tend to do better at replenishing energy lost to exhaust gas flow. E-core topologies have an advantage of lower moving armature mass for a design with the same footprint and the same latching force. Further pole shaping issues will be discussed below.
The second related design issue concerns a quick transition from armature deceleration to magnetic latching. If the magnetic force attracting an armature builds smoothly to a maximum value at latching, the implication is that the magnetic force will approach a balance with opposing mechanical spring force, causing deceleration of the armature to slow as closure is approached along a smooth, gradual curve. For closing an automotive valve xe2x80x9ccrisplyxe2x80x9d without a lengthy and gradual approach, the magnetic pull-in force should ideally be quite low as the valve approaches closure, so that the spring force will be almost unopposed in decelerating the valve on approach to the valve seat. The problem is that if the armature and valve close while magnetic force is low and spring-driven deceleration is high, then the valve will start to re-open rapidly, before the magnetic flux can be built up to a sufficient level to latch the valve shut. Even though little current and little power are required for latching, the problem is a limited slew rate of magnetic flux linkage, a rate that is limited by the peak available coil drive voltage. Looking at it differently, one can say that the peak available coil drive voltage sets an upper limit on the controllable third derivative of valve position, which is the rate of change of acceleration. Smooth but prompt landing requires a large swing of the third derivative of motion. In this regard, the latching voltage pulse illustrated on trace 430 between time markers 495 and 496 is a difficult issuexe2x80x94one would like a much faster armature approach with a much lower flux, ending in a much bigger latching pulse. In the simulation for FIG. 4, magnetic flux trace 450 is maintained at or near its saturation limit for an extended period, providing pull to replenish valve and armature mechanical energy lost to exhaust gas forces, and thus keeping the valve moving toward full-open with solenoid closure, so that the motion does not fall short of latching. Since the electromagnetic pulling force is high even between time markers 494 and 495, the decelerating spring force is largely neutralized and the braking deceleration of trace 420 is low, giving a slow approach to closure. In a typical case where there is less energy loss to make up, particularly in an intake valve or any valve moving from open to closed position (i.e. the reverse direction of the valve-open stroke illustrated in FIG. 4), the magnetic flux could be quite low over the range of armature positions (trace 420) encountered between time markers 493 and 495. While this would give a faster approach with greater deceleration, the latching pulse on trace 430 initiated at time marker 495 would have to start much earlier, in order to pull flux linkage trace 450 up to the needed latching level starting from a much lower level during mid-trajectory. Thus, the quickness gained from less gas flow energy loss and a higher kinetic energy of the valve and armature crossing the middle range of motion, is mostly sacrificed during a more extended latching pulse as the latching transition marked by 495 is moved earlier in time.
To illustrate the magnitude of the quickness problem just described, a sinusoidal motion at the natural period of the spring and effective moving mass of the valve, armature, and spring simulated in FIG. 4 would give a half-period transition in 3.3 milliseconds. The actual transition time from initial motion just before time marker 491 to the first small touch-down bounce at time marker 496 is roughly 6.3 milliseconds, or about 1.9 times longer than a xe2x80x9cnominalxe2x80x9d transition time computed from the spring and mass alone. This 1.9-times multiplier becomes somewhat smaller where there is less energy loss to make up, but the figure remains well above 1.5 for most practical actuator designs.
The winding impedance matching problem previewed above is now seen in context. Drive voltage and current are both quite high on approach to time marker 493, as energy losses are replenished under low efficiency conditions. Given a limited power supply voltage, e.g., 42 volts DC, and given the expense of transistor circuitry to handle more than 40 or 50 amperes efficiently at such a low voltage, the designer is constrained to design a yoke winding with enough turns to give a resistive impedance of at least a few tenths of one ohm. For very rapid flux slewing with a small magnetic gap, where solenoid inductance is high and is the current-limiting factor in the short term, one would like a winding with very few turns, winding resistance not being an issue. When xe2x80x9cvery fewxe2x80x9d is quantified, the outcome is that one wants fewer turns for quick latching than one wants for high pull-in power. When the turns count is constrained to achieve high net magnetic pull-in power, valve closure that is both prompt and low impact becomes difficult or impossible to achieve. Too few volts and too many winding turns implies a cap on the third derivative of armature motion on approach to closure, and soft landing is achieved only by a slow and xe2x80x9clazyxe2x80x9d motion. Quicker landing is achieved only with impact. If the bounce resulting from impact is not sufficiently damped, the armature will bounce so far open that magnetic forces cannot prevent the solenoid from failing to latch.
Various electronic schemes can permit high peak voltages at low currents to achieve quicker latching, but at a cost. The product of peak volts times peak amps from an amplifier entails a cost, even if the peak instantaneous wattage is substantially less than the product of a peak voltage occurring at a different time than the peak amperage. By conventional means, therefore, it can be difficult to design a valve solenoid that makes up substantial energy losses, especially in an exhaust valve, and simultaneously achieves very prompt latching.
An object of the invention is improvement of passive electromagnetic characteristics in closing and magnetically latching solenoids, contributing to overall efficiency, stability, and control in several areas: for pulling magnetically across a large gap to maintain kinetic energy and pull an armature to a closed latching position against a powerful spring; for electronically controlled soft landing; for a rapid increase of latching force created passively as a function of the mechanical closure of armature with yoke; for a rapid decrease of latching force created passively as a function of the mechanical opening of armature with yoke; and for reduction of bounce following closure of armature with yoke. A related object is to improve the efficiency with which a solenoid pulls across a large gap by enlarging the ferromagnetic attraction area between armature and yoke. A further related object is to provide one or more recesses in the attraction area, so that on closing or opening transitions, the magnetic flux through the solenoid abruptly shifts away from or toward the recesses and toward or away from a reduced poleface area achieving final mating contact, resulting in an abrupt passive increase or decrease in magnetic force on approach to or withdrawal from closed contact. As a means of achieving the above-described objects of enlarging the attraction area at a distance and causing flux to shift to a reduced mating poleface area for final mating contact, while at the same time minimizing the inertial mass of the moving armature, it is an optional object to recess some poleface area to prevent mating and to concentrate the reduced mating poleface area toward the middle gap or gaps, respectively, between the two legs of a U-core yoke or the three legs of an E-core yoke, and to reduce the armature thickness outside this reduced mating area, where less magnetic flux capacity is required. As an alternative means for enlarging attraction area at a distance and causing flux shift to a reduced mating poleface area, it is an optional object to create a multiplicity of narrow or small recesses distributed across the poleface area. Finally to achieve passive electromagnetic damping of magnetic closure or opening with flux shift as described above, it is an object to provide intentional eddy current conduction paths, responsive to the flux shift upon solenoid closure or opening by generating induced electric currents and dissipating the energy of those currents resistively. Optionally, a related object is to create the intentional eddy current conduction paths in the topology of one or more figure-8 windings, whose induced current response is low for a change in drive winding flux linkage at a large armature-yoke gap, but whose induced current response for a fixed drive winding flux linkage is high when flux paths shift laterally upon mechanical closure or opening of the armature-yoke gap. A further optional object related to passive electromagnetic damping is to control the presence or phase response or direction of that damping by insertion of an electrical impedance element or controllable impedance element, such as a capacitor or diode or transistor, in the eddy current conduction path.
In the prior art, magnetic force F is commonly modeled in terms of the square of magnetic flux "PHgr" that links the solenoid windings, multiplied by a variable function f of position X, according to Eq. 1:
F=xe2x88x92"PHgr"2xc2x7f(X)xe2x80x83xe2x80x831]
The negative sign indicates a force of attraction, toward decreasing X, while X is commonly defined so that X=0 corresponds to mating of magnetic pole faces. In many treatments, the function f(X) is approximated as a constant coefficient, which is a fair approximation for some solenoids with long armatures, but which is a poor approximation for E-core, U-core, and pot core topologies with short armatures, as commonly applied to internal combustion intake and exhaust valve actuators, for example. Where the variability of f(X) is considered in the prior art, it is taken as a given, not easily altered within the geometric constraints of a problem. In control design, therefore, the focus has been on electrically altering flux "PHgr" adaptively during the travel of the armature, to control motion, correct for gas-flow-related gains or losses of energy in transit, and achieve soft landing and latching. An engineering problem associated with this focus is the difficulty of varying flux "PHgr" rapidly for soft landing and latching while maintaining efficiency for transferring substantial amounts of energy to or from the armature, to compensate for gas flow effects.
The present invention describes a method and device design, in multiple embodiment variations, for altering poleface geometry in order to tailor the attraction function f(X) to meet control needs, especially for combinations of improved electromechanical efficiency of energy transfer and of creating variations in f(X), on approach to landing and latching, that permit quick landing and latching within slewing and saturation limits associated with "PHgr". The invention utilizes an expanded poleface area for efficient magnetic attraction to achieve energy transfer at a distance, combined with an abrupt increase in f(X) as gap X approaches zero. This abrupt increase in f(X) causes a large third derivative of armature motion which, on landing, gives a rapid transition from deceleration to a stronger attractive force for latching. On armature release, the same abrupt change in f(X) causes a rapid release from a latching condition with transition to high acceleration.
The physical mechanism behind the abrupt increase in f(X) as X approaches zero is understood in relation to Eq. 2, which expresses the normal force Fn attracting a pair of flat parallel surfaces of ideal ferromagnetic conductor, mating over a net area A and conducting net flux "PHgr" across a very small gap over the mating area (i.e., expressing force in the limit as gap Xxe2x86x920):
Fn=xe2x88x92"PHgr"2/(2xcexcoA)xe2x80x83xe2x80x832]
Note that force F in Eq. 1 is a net axial force exerted on an armature, whereas the normal force Fn may be for a fraction of an armature poleface area whose normal vector is not parallel to the axis of armature motion. Also, Eq. 2 is accurate only in the limit as the gap between the parallel surfaces approaches zero. It will nevertheless be recognized from Eqs. 1 and 2 that the force scaling function f(X) will be increased if the effective flux-carrying poleface area A can be caused to decrease with decreasing axial gap X.
Eqs. 1 and 2 are instructive where magnetic flux is restricted in some way, either by saturation or by slew rate. A different way of writing these expressions focuses attention on power consumption, a useful perspective when flux is not a limitation. Eq. 3 is closely related to Eq. 1:
F=xe2x88x92I2xc2x7g(X)xe2x80x83xe2x80x833]
Since resistive power dissipation is given by Pwr=I2xc2x7R for resistance R, Eq. 3 indicates power efficiency in producing a magnetic force. Power dissipation is inversely related to g(X) for a given force F. Eq. 2 indicated that the flux-factor f(X) is inversely related to area. Eq. 4 indicates that the current-factor g(X) is proportionally related to area:
Fn=xe2x88x92I2n2xcexcoA/(2X2)xe2x80x83xe2x80x834]
Like Eq. 2, Eq. 4 is accurate only for small values of gap X, but it indicates important trends. Solenoid power consumption varies with the square of the ampere-turn product, I2n2, which is intimately related to power for a given winding window. While a change in wire size affects the number xe2x80x9cnxe2x80x9d of windings that will fit in a window, when one accounts for the resistance per turn of the wire that will fit the window, one finds that power varies fairly accurately with I2n2, regardless of the wire gauge used. Thus, Eq. 4 indicates that an increase in effective magnetic attraction area A improves power efficiency, whereas Eq. 2 shows that a decrease in A increases force when flux "PHgr" is the limiting factor. These properties are reflected in the functions f(X) and g(X) of Eqs. 1 and 3, such that f(X) and g(X) are oppositely related to magnetic attraction area. For obtaining a maximum latching force, the mating poleface area should be reduced to the minimum size that will conduct all the flux that can be carried by the remainder of the magnetic circuit. Magnetic saturation places a practical upper bound on "PHgr", while limitations on voltage applied to a solenoid drive winding set a practical slew rate limit on "PHgr". Thus, in the last few hundred microseconds before a solenoid closes, flux "PHgr" cannot change significantly, so Eqs. 1 and 2, focusing on a fixed flux, best describe the force curve realizable in a very short time frame on landing and latching. When power consumption rather than flux is a constraining situation, Eqs. 3 and 4 focus best on the situation.
The multiple demands on a quick latching solenoid design are summarized as follows. First, the effective attraction area xe2x80x9cAxe2x80x9d should be kept as large as possible for large gaps X to provide efficient pull at distances where current limitations make it difficult or impossible to reach saturation limits. Second, the effective area A should be reduced abruptly as X approaches zero, in order to achieve a quick force increase at constant flux and switch from deceleration to latching. Third, the reduced mating poleface area at latching should be no larger than necessary to carry the flux of the remainder of the magnetic circuit at the threshold of saturation.
In solenoids with simple flat polefaces, as in FIGS. 1, 2, and 3, the magnetic field between widely spaced polefaces actually spreads out to cover an effective area larger than the poleface surfaces, and this effective area shrinks as the polefaces close and X approaches zero. Thus, f(X) in Eq. 1 is almost always an increasing function of decreasing X approaching zero. This increasing trend in f(X) as X approaches 0 is made even greater by the progressive reduction of stray flux linking the drive winding of a solenoid but shorting between parts of the yoke without bridging across to the armature. Modifying this typical trend, the behavior of f(X) can be substantially altered by designing for a lateral redistribution of magnetic flux across pole faces approaching mating closure, as portions of the pole faces are slightly recessed so that more prominent portions of the pole faces mate and draw flux away from the recessed or non-mating portions. Narrow grooves can be used in place of broader areas of shallow recess to redistribute flux away from the grooves as X approaches zero.
In the above description, flux shift of two sorts occurs. Solenoids lacking flux shift features, as illustrated in FIGS. 1, 2, and 3, have axial-facing or nearly axial-facing mating poleface area, and laterally-facing side area. The poleface area on the yoke includes regions of differing magnetic polarity, some area characterized by magnetic-north polarity and other area characterized by magnetic-south polarity, depending on the current rotation sense in the drive winding of the yoke. As a gap opens between mating axial-facing areas with armature motion away from the yoke, flux begins to flow through side areas of the yoke, with some of this flux leaking across from one side of the yoke to another, e.g. from a north-polarity to a south-polarity region, without entering and leaving the armature. This flux shift from facing area to side area reduces the axial armature force at a given net flux linkage from the yoke winding or windings. Solenoids including intentional flux shift features include, in addition to side area and axial-facing mating poleface area, additional axial-facing non-mating poleface area, which is recessed in some way. This non-mating area is typically shaped to add to the effective attraction area at a large gap from armature to yoke, resulting in more flux bridging from the yoke to the armature for a given magnetomotive force of winding ampere-turns. The force per ampere-turn is thus higher, but the flux is also comparatively high. At short ranges, where saturation commonly limits the total flux that can go to the armature, a concentration of the available flux into a smaller mating poleface area results in a higher latching force. The abruptness of the transition from efficient pull-in to high latching force at saturation can be controlled, making it abrupt for prompt landing, or making the transition more gentle if continuous servo control is sought down to very small armature-yoke gaps. A more gradual flux shift reduces destabilizing magnetic force gradients with respect to changing armature position, thus easing the demand on servo control. One might say that the recessed area is intentional by design, according to principles taught herein, and not an accidental design artifact, if less than 90% of the attracting poleface area is mating, with the remainder being axially facing but non-mating and recessed by more than an artifact of fabrication errors. For a solenoid characterized by xe2x80x9crapidxe2x80x9d flux shift, a small axial gap opening, say 10% of the maximum axial armature travel, will produce a disproportionately large reduction in force at constant flux linkage through the drive winding, e.g., a force reduction of 20% or more. Without design for intentionally augmented flux shift, a 10% armature travel would typically produce a force reduction significantly less than 20%. These figures are not intended as limiting, but are given by way of example of magnitudes to be expected in common engineering designs.
Another way to achieve flux redistribution in an E-core topology is to provide lateral flux paths across gaps on the sides of the armature. The lateral flux does not generate much axial force, but the addition of low-reluctance lateral paths to side plates, bridging between the outside legs of the E-core and the sides of the armature, promotes greater total flux conduction for the same magnetomotive force. The effect is to increase the pull between the armature and the center post of the E-core for a given number of ampere turns, thus increasing the efficiency of pull at a distance. As the armature approaches axial closure with the E-core, the axially facing surfaces at the ends of the armature close with axially facing surfaces of the outer legs of the E-core, drawing flux away from the lateral gaps and creating an abrupt or gentle increase in axial force at constant flux linkage in the winding, depending on the geometric proportions. By combinations of lateral or oblique flux conduction paths, recessed poleface surfaces, and surfaces that mate, working with E-core, U-core, pot core, and other topologies, one can promote efficient pull-in across large gaps and control the increasing curve of force with decreasing poleface gap. This force control can include very steep increases in force with decreasing gap at small gaps, or more moderate rates of increase in force, depending on design tradeoffs between stable servo control and quick latching in time frames too short for effective servo control. By way of numerical example, but without limitation of the invention, when the armature in a typical design with ferromagnetic side plates is midway in its travel between the dual E-cores, a substantial fraction of flux entering the armature from the center pole of an electrically-energized E-core will travel laterally out of the armature and into the side plates, e.g., 50% or more into the side plates. (This is true only if one of the two E-cores is energized by a winding current. If both E-cores carried current, e.g. equal currents, then flux from one E-core center pole to the armature might pass through the armature and be drawn to the opposite E-core center pole, rather than flowing laterally to the side plates.) If the fractional flux into the side plates is much less than 50%, one might say that the side plates could be providing comparatively little advantage for efficient pull-in from a distance. When the armature travels to latch to the energized E-core, most of the flux that was traveling laterally into the side plates ought to shift to an axial path into axially facing surfaces of the outer posts of the E-corexe2x80x94e.g., with 80% or more of the flux shifted to the axial path. As with earlier figures, these figures are provided not by way of limitation, but by way of example of magnitudes associated with useful flux shift, for providing a significant improvement in the combination of maximum holding force at saturation, and efficiency at generating force across a large gap with reduced resistive power loss.
In either E-core or U-core or pot core solenoid topologies, the flux-carrying material in an armature must carry maximum flux across flux gaps, e.g. between the legs of an E-core or U-core, or in the radial gap of a pot core topology. Less flux is carried elsewhere, far from the flux gaps. The amount of flux-carrying material is varied across the width (or radius) of the armature to roughly the minimum, at each point, needed to avoid local saturation. This variation in flux-carrying material can be accomplished by stepping or tapering the portions of the armature that are away from flux gaps, or by creating cavities inside parts of the armature away from flux gaps. The armature mass is thus minimized. By further concentrating recessed areas away from flux gaps, the armature mass can be minimized in expanded, recessed areas. Thus, one achieves efficient attraction at moderate flux levels for large X, utilizing laterally extended and axially recessed poleface areas; one achieves an abrupt increase of force on approach to landing with a redistribution of flux inward to the thick, high-flux-capacity portions of the armature bridging gaps; and one maintains a moderately low armature mass.
Further modifications of the design provide intentional eddy current paths, associated geometrically with the areas of lateral flux redistribution for the abrupt increase in f(X), so that energy is dissipated during the flux redistribution, and damping is introduced into the force transition from deceleration to latching. Through modifications of f(X) and of damping associated with rapid changes in f(X) with changes in X, a designer applying the teachings of the current invention can create a target window, at a given small Xo, described in terms of the control space variables of velocity dX/dt and flux "PHgr". If, upon reaching Xo with a negative dX/dt (i.e. with the gap closing), the control space position (dX/dt, "PHgr" lies within the target window, then without further application of coil drive voltage, the armature will land without excessive impact or bounce and will become latched. The idea is that for an Xo so small and so close to the landing point that further active course correction is ineffective, the poleface design leading to f(X) and damping associated with dynamic changes in f(X) will cause the armature to achieve a successful landing and latching. The electromagnetic design of the solenoid substantially takes over the control process from the electronic servo controller in order to land the armature quickly. This quick landing is achieved successfully if the servo controller can hit a finite landing target window. The processes that occur on landing in such a design are so quick, and involve such large redistributions of magnetic flux, that they would be difficult to achieve through active electronic control in the absence of an improved passive electromagnetic design, as described in more detail below.